ML20023B693

Rev 1 to Design Rept for Recirculation Sys & RHR Sys Weld Overlay Repairs & Flaw Evaluation at Ei Hatch Nuclear Power Plant,Unit 1.

ML20023B693
Person / Time
Site:	Hatch
Issue date:	03/04/1983
From:	Charnley J, Riccardella P NUTECH ENGINEERS, INC.
To:
Shared Package
ML20023B692	List: ML20023B690 ML20023B693
References
GPC-04-104, GPC-04-104-R01, GPC-4-104, GPC-4-104-R1, GPC004.0104, GPC4.0104, TAC-49156, NUDOCS 8305060055
Download: ML20023B693 (100)

Text

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GPC-04-104 Revision 1 March 1983 GPC004.0104 DESIGN REPORT FOR RECIRCULATION SYSTEM AND RESIDUAL HEAT REMOVAL SYSTEM WELD OVERLAY REPAIRS AND FLAW EVALUATION AT E. I. HATCH NUCLEAR POWER PLANT UNIT 1 Prepared for Georgia Power Company Prepared by NUTECH Engineers, Inc.

San Jose, California Prepared by: Reviewed by:

W Ere J. E. Charnley, P.E. P. E. Reeves j Project Engineer Project Quality Assurance Engineer Q '

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l Approv;ed/b y: g Issued by:

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7 Senior Director Project Manager Date: f/A,ud 4,. /983 8305060055 030428 PDR ADOCK 05000321 nutech.

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REVISION CONTROL SHEET Design Report for Recirculation System TITLE; and Residual Heat Removal System Weld REPORT NUMBER: GPC-04-104 Overlay Pepairs and Flaw Evaluation Revision 1 at Hatch 1 J. E. Charnley / Principal Engineer _,

P. C. Riccardella/ Senior Director Y. S. Wu/ Consultant I [

K. W. Benting/Cor.sultant I Mh J. R. Taylor / Consultant II h/'7" T. Lem/Ccasultant I [

R. D. Carignan/ Principal Encineer H. L. Gustin / Engineer NAME / TITLE INITIALS PA ACWRACY NERIA REMARKS PAGE(5) REV SY / CATE CHECX BJ / DATE CHECK 8pDATE ii 111 o

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REVISION CONTROL SHEET Design Report for (CONTINUATION)

TITLE: Recirculation System and Residual Heat REPORT NUMBER: GPC-04-104 Removal System Weld Overlay Repairs Revision 1 and Flaw Evaluation at Hatch 1 PREPARED ACCURACY CRITERIA

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l REVISION CONTROL SHEET Cesign Report for (CONTINUATION)

TITLE: Recirculation System and Residual Heat REPORT NUMBER: GPC-04-10 4 Re:nov.tl System Weld Overlay Repairs Revision 1 and F2aw Evaluation at Hatch 1 PREPAAED ACCUP.ACY CRITERIA PAGEIS) AEV REMARC SY/CATE CHECX,8Y / CATE CHE):3, pry / day 71 72 0 h% INS lE3 ({0L Ik8$ [!NbIfG 73 74 75 76 77 78 79 80 81 82 83 84 85 ' U f

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CERTIFICATION BY REGISTERED PROFESSIONAL ENGINEER I hereby certify that this document and the calculations contained herein were prepared under my direct supervision, reviewed by me, and to the best of my knowledge are correct and complete. I am a duly Registered Professional Engineer under the laws of the State of California and am competent to review this document.

Certified by:

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GPC-04-104 v Revision 1 nut.ech

TABLE OF CONTENTS Page LIST OF TABLES vii LIST OF FIGURES viii

1.0 INTRODUCTION

1 2.0 REPAIR DESCRIPTION 4 3.0 EVALUATION CRITERIA 8 3.1 Weld Overlay Evaluation 8 3.1.1 Strength Evaluation 9 3.1.2 Fatigue Evaluation 9 3.1.3 Crack Growth Evaluation 10 3.2 Flaw Evaluation 11 4.0 LOADS 12 4.1 Mechanical and Internal Pressure Loads 12 4.2 Thermal Loads 12 5.0 EVALUATION METHODS AND RESULTS 14 5.1 End Cap Evaluation 14 5.1.1 Code Stress Analysis 15 5.1.2 Fracture Mechanics Evaluation 17 5.2 Elbow Evaluation 24 5.2.1 Code Stress Analysis 24 5.2.2 Fracture Mechanics Evaluation 26 5.3 Pipe-to-Pipe Evaluation 31 5.3.1 Code Stress Analysis 32 5.3.2 Fracture Mechanics Evaluation 33 L

5.4 Sweepolet Evaluation 37 5.4.1 Code Stress Analysis 38 5.4.2 Fracture Mechanics Evaluation 39 5.5 Effect on Recirculation and RHR Systems 43 GPC-04-104 vi Revision 1 nutp_qh

TABLE OF CONTENTS (Continued)

Page 6.0 LEAK-BEFORE-BREAK 69 6.1 Net Section Collapse 69 6.2 Tearing Modulus Analysis 70 6.3 Leak Versus Break Flaw Configuration 71 6.4 Axial Cracks 72 6.5 Multiple Cracks 73 6.6 Crack Detection Capability 73 6.7 Non-Destructive Examination 74 6.8 Leakage Detection 75 6.9 Historical Experience 76 7.0

SUMMARY

AND CONCLUSIONS 82

8.0 REFERENCES

84 9.0 ENGINEERING CHANGES 87 i

GPC-04-104 vii Revision 1 nutp_qh

LIST OF TABLES Number Title Page 5.1 Thermal Stress Results 45 5.2 End Cap Code Stress Results 46 5.3 Elbow Code Stress Results 47 5.4 Pipe-to-Pipe Code Stress Results 48 6.1 Effect of Pipe Size on the Ratio 77 of the Crack Length for 5 GPM Leak Rate and the Critical Crack Length (Assumed Stress o= (Sm)/2) r i

GPC-04-104 viii Revision 1 nut.e_qh

LIST OF FIGURES Number Title Page 1.1 Conceptual Drawing of Recirculation 3 and RHR Systems 2.1 Schematic of End Cap Weld Overlay 5 2.2 Schematic of Elbow-to-Pipe Weld Overlay 6 2.3 Schematic of Pipe-to-Pipe Weld Ove'rlay 7 5.1 End Cap Finite Element Model 49 5.2 Weld overlay Thermal Model 50 5.3 Thermal Transients 51 5.4 Axial Crack Growth Residual Stress 52 5.5 Typical IGSCC Crack Growth Data 53 (Weld Sensitized 304SS in BWR Environment) 5.6 End Cap Axial Crack Growth 54 5.7 End Cap Tearing Modulus 55 5.8 Elbow Finite Element Model 56 5.9 Circumferential Crack Growth Residual Stress 57 5.10 Elbow Circumferential Crack Growth 58 l

5.11 Elbow Tearing Modulus 59 5.12 Pipe-to-Pipe Finite Element Model 60 5.13 Pipe-to-Pipe Axial Crack Growth 61 5.14 Pipe-to-Pipe Tearing Modulus 62 7 5.15 Sweepolet Crack Geometry 63 5.16 Sweepolet Finite Element Model (Outside) 64 5.17 Sweepolet Finite Element Model (Inside) 65 5.18 Sweepolet Crack Growth 66 GFC-04-104 ix Revision 1 nutggb

LIST OF FIGURES (Continued)

Number Title Page 5.19 Sweepolet Tearing Modulus 67 5.20 Piping Model 68 6.1 Typical Result of Net Section Collapse 78 Analysis of Cracked Stainless Steel Pipe 6.2 Stability Analysis for BWR Recirculation 79 System (Stainless Steel) 6.3 Summary of Leak-Before-Break Assessment 80 of BWR Recirculation System 6.4 Typical Pipe Crack Failure Locus for Combined 81 Through-Wall Plus 360* Part-Through Crack 9.1 Sketch of BR-13 Overlay Toe Crack 88 9.2 Sketch of Revised SR-13 Overlay 89 9.3 Revised Pipe-to-Pipe Finite Element Model 90 f

GPC-04-104 x Revision 1 nutp_qh

1.0 INTRODUCTION

This report summarizes evaluations performed by NUTECH to assess weld overlay repairs and unrepaired flaws in the Recirculation and Residual Heat Removal (RHR)

Systems at Georgia Power Company's Edwin I Hatch Nuclear Plant Unit 1. Weld overlay repairs have been applied to l address leakage and additional ultrasonic (UT) examination results believed to be indicative of intergranular stress corrosion cracking (IGSCC) in the vicinity of the welds. The purpose of each overlay is to arrest any further propagation of the cracking, and to restore original design safety raargins to the weld.

The unrepaired welds which had UT examination indications have been shown by analysis to continue to have the original design safety margins.

The required design life of each weld overlay repair is at least two fuel cycles. The amount that the actual design life exceeds two fuel cycles will be established I

by a combination of. future analysis and testing.

f Leakage was observed during overlay welding adjacent to one end cap-to-manifold weld and in addition, crack indications have been detected adjacent to three end GPC-04-104 1 Revision 1 nute_qh

cap-to-manifold welds, one elbow-to-pipe weld, one pipe-to-pipe weld and one sweepolet-to-manifold weld, All of these welds except the sweepolet-to-manifold weld were repaired with weld overlay designs evaluated in this report. The analysis of the unrepaired sweepolet weld is also contained herein.

Figure 1.1 shews the end caps, elbow, sweepolet and pipe-to-pipe welds in relation to the reactor pressure l

vessel and other portions of the recirculation and RHR systems. All of the existing material is type 304 stainless steel.

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GPC-04-104 2 Revision 1 nut _ech

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,=e= CONCEPTUAL DRAWING OF RECIRCULATION AND RilR SYSTEMS

2.0 REPAIR DESCRIPTION The through-wall cracks and other indications around and to both sides of the existing end cap, pipe-to-pipe and elbow weld heat-af fected zones have been repaired by establishing additional " cast-in-place" pipe wall thickness from weld metal deposited 360 degrees around and to either side of the existing weld, as shown in Figures 2.1, 2.2, and 2.3. The weld deposited band over the cracks will provide wall thickness equal to that required to provide the original design safety margins. In addition, the weld metal deposition will produce a favorable compressive residual stress pattern. The deposited weld metal will be type 308L, which is resistant to propagation of IGSCC cracks.

GPC-04-104 4 Revision 1 nutech

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3.0 EVALUATION CRITERIA 3.1 Weld Overlay Evaluation This section describes the criteria that are applied in this report to evaluate the acceptability of the weld overlay repairs described in Section 2.0. Because of the nature of these repairs, the geometric configuration is not directly covered by Section III of the ASME Boiler and Pressure Vessel Code, which is intended for new construction. However, materials, fabrication procedures, and Quality Assurance requirements are in 1

accordance with applicable sections of this Construction Code, and the intent of the design criteria described below is to demonstrate equivalent margins of safety for strength and fatigue considerations as provided in the ASME Section III Design Rules. In addition, because of the IGSCC conditions that led to the need for repairs, IGSCC resistant materials have been selected for the weld overlay repairs. As a further means of ensuring structural adequacy, criteria are also provided below for fracture mechanics evaluation of the repairs.

GPC-04-104 8 Revision 1 nut _ech

3.1.1 Strength Evaluation Adequacy of the strength of the weld overlay repairs with respect to applied mechanical loads is demonstrated with the following criteria:

1. An ASME Boiler and Pressure Vessel Code Section III, Class 1 (Reference 1) analysis of the weld overlay repairs was performed.

2. The ultimate load capacity of the repairs was calculated with a tearing modulus analysis. The l ratio between failure load and applied loads was required to be greater than that required by Reference 1.

3.1.2 Fatigue Evaluation The stress values obtained from the above strength evaluation were combined with thermal and other secondary stress conditions to demonstrate adequate fatigue resistance for the design life of each repair.

The criteria for fatigue evaluation include:

GPC-04-104 9 Revision 1 nutegj)

1. The maximum range of primary plus secondary stress was compared to the secondary stress limits of Reference 1.

2. The peak alternating stress intensity, including all primary and secondary stress terms, and a fatigue strength reductic.1 factor of 5.0 to account for the existing crack, was evaluated using conventional fatigue analysis techniques. The total fatigue usage factor, defined as the sum of the ratios of applied number of cycles to allowable number of cycles at each stress level, must be less than 1.0 for the design life of each repair.

Allowable number of cycles was determined from the stainless steel fatigue curve of Reference 1.

3.1.3 Crack Growth Evaluation Crack growth due to both fatigue (cyclic stress) and IGSCC (steady state stress) was calculated. The allowable crack depth was established based on net section limit load for each cracked and repaired weld (Reference 2). The design life of each repair was GPC-04-104 10 Revision 1

established as the minimum of either the predicted time for the observed crack to grow to the allowable crack depth or five years.

3.2 Flaw Evaluation Crack growth due to both fatigue (cyclic stress) and IGSCC (steady state stress) was calculated. The allowable crack depth was established based on the net section limit load for the cracked sweepolet weld (Reference 2).

The life of the sweepolet weld with the observed flaw was established as the time for the flaw to grow to the allowable crack depth.

l GPC-04-104 11 Revision 1 nut _ech

4.0 LOADS The loads considered in the evaluation of the sweepolet, end cap, pipe-to-pipe and elbow welds consist of mechanical loads, internal pressure, differential thermal expansion loads, and welding residual stresses. The mechanical loads and internal pressures used in the analysis are described in Section 4.1, and an explanation of the thermal transient conditions which cause differential thermal expansion loads is presented in Section 4.2. Welding residual stresses are considered in the crack growth analyses and are described in Sections 5.1.2.2 and 5.2.2.2.

4.1 Mechanical and Internal Pressure Loads The design pressure of 1325 psi for the Recirculation and RHR Systems was obtained from Reference 3. The dead weight and seismic loads applied to each weld were obtained from Reference 4.

4.2 Thermal Loads i

The thermal expansion loads for each weld were also obtained from Reference 4 and applied to the weld GPC-04-104 12 Revision 1 nut _ech

overlay repairs. Reference 3 defines several types of transients that the Recirculation and RHR Systems are designed for. These transients were conservatively grouped into three composite transients. The firet composite transient is a startup/ shutdown transient with a heatup or cool down rate of 100*F per hour. The second composite transient consists of a 50 F step temperature with no change in pressure. The third composite transient is an emergency event with a 416*F step temperature change and a pressure change of 1325 psi. In the five year overlay design life, there are 38 startup/ shutdown cycles, 25 small temperature change cycles, and 1 emergency cycle.

1 l

GPC-04-104 13 Revision 1 nutggh

5.0 EVALUATION METHODS AND RESULTS, The evaluation of the weld overlay repairs and the unrepaired sweepolet weld consists of a code stress analysis per Section III (Reference 1) and a fracture mechanics evaluation per Section XI (Reference 5).

5.1 End Cap Evaluation The four end cap welds with ultrasonic indications are:

1) 1B31-lRC-22-AM-1 with maximum depth of 63% of wall

2) 1B31-lRC-22-AM-4 with maximum depth of 72% of wall

3) 1B31-lRC-22-BM-1 with maximum depth of 64% of wall

4) 1B31-lRC-22-BM-4 with maximum depth of 67% of wall All indications are axial, with a length of approxi-mately 1/2".

During the installation of the weld overlays, porosity l

appeared on the exterior surface of the AM-4, BM-4, and 1

l BM-1 end cap welds. The locations of the porosity were recorded and the weld overlays completed. The porosity was then repaired by grinding back to base metal and filling the cavity with shielded metal arc welding.

GPC-04-104 14 Revision 1 nut.e_qh

5.1.1 Code. Stress Analysis The weld overlaid regions were assumed to be

- axisymmetric. That is, a 75% through-wall axial crack s

was conserve.tively assumed to be 360 degrees around the pipe and 1/2 inch long on the end cap side of the weld. It is judged that the assumed crack geometry conservatively envelopes all observed cracks in the end cap welds. In cddition, all analyses were conserva-tively performed using a weld overlay thickness of 0.25 inch which,is'10 percent smaller than the actual minimum average thickness of 0.275 inch. A finite element model of the cracked and weld overlaid region was developed using the ANSYS (Reference 6) computer program. Figure 5.1 shows the model.

Based on Reference 4, the applied thermal, weight and seismic moments on these welds are negligible. The s

pressure stress for a design pressure of 1325 psi was calculated with this model. The weld overlay thermal model was also taken to be axisymmetrical (Figure 5.2). The exterior boundary was assumed to be insulated. The temperature distribution in the weld overlay subject to the thermal transients defined in Section 4.2 can be readily calculated using Charts 16 GPC-04-104 15

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9 and 23 of Reference 7. The maximum through-wall 4

s temperature difference was determined to be less than 2*F for the normal startup cycle, 40*F for the small temperature cycle and 329'F for the emergency transient.

The maximum thermal stress for use in the fatigue crack growth analysis was calculated as follows: (Reference 1)

Ea aT Ea aT 7 2

+

s = 2 ( 1- v) l- v Where:

6 E = 28.3 x 10 poi (Young's Modulus) a = 9.11 x 10-6 op-l (Coefficient of Thermal Expansion)

Er = Equivalent Linear Temperature Difference dr 2 = Peak Temperature Difference The values of dr y , dr , and s are given in Table 5.1 for 2

all three thermal transients.

i, The results of a code stress analysis per Reference 1 are given in Table 5.2. The allowable stress values for i

GPC-04-104 16 Revision 1 nutggb

t Reference 1 are also given. The weld overlay repair i satisfies the Reference 1 requirements.

A conservative fatigue analysis per Reference 1 was performed. In addition to the stress intensification factors required per Reference 1, an additional fatigue strength reduction factor of 5.0 was applied due to the crack. The fatigue usage factor was then calculated assuming 38 startups, 25 small temperature change cycles and one emergency cycle every five years. The results are summarized in Table 5.2.

5.1.2 Fracture Mechanics Evaluation Three types of f racture mechanics evaluations were performed. The allowable crack depth was calculated based on Reference 2. Crack growth due to both fatigue and IGSCC was calculated using the NUTECH computer program NUTCRAK (Reference 8) with material constants and methodology from References 9 and 10. Finally, the ,

i ultimate margin to failure for a crack assumed to  !

propagate all the way through the original pipe material to the weld overlay was calculated per References 11 and 12.

GPC-04-104 17 Revision 1 nutggb

5.1.2.1 Allowable Crack Depth The allowable depth for a 1/2 inch long axial crack was determined using Reference 2. The dimensions of the repaired pipe were used. Thus, the ratio of applied primary stress to Code allowable stress (Sm) was calculated in the follcwing manner:

Stress Ratio = PR/t g

m P = 1325 psi (Design Pressure)

R = 11.20 inches (Outside Radius of Pipe) t = 1.24 inches (Nominal Pipe Thickness)

S, = 16,800 psi (Reference 1)

Substitution yields:

Stress Ratio = .71 The nondimensional crack length (3) was calculatad in the following manner:

GPC-04-104 18 Revision 1 nutech

L g

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(rt)1/2 t

g

= 1/2 inch r = 10.58 inches (mean radius of pipe) t = 1.24 inches Substitution yields:

Nondimensional Length = .14 Thus per Table IWB-3642-1 of Reference 2, the allowable crack depth is 75 percent of the wall thickness.

5.1.2.2 Crack Growth The existing cracks could grow due to both fatigue and stress corrosion. Fatigue crack growth due to the three types of thermal transients defined in Section 4.2 was calculated using the material properties from Reference 9. The fatigue cycles considered are shown in Figure 5.3. The fatigue crack growth for 5 years was calculated to be less than 0.01 inch.

GPC-04-104 19 Revision 1 nutagh

IGSCC growth depends on the total steady state stress.

The major contributor to the steady state stress is the weld residual stress. The residual stress due to the original butt weld was conservatively chosen to be a worst case with through-wall bending stress of 30,000 psi with tension on the inside surface. The weld residual stress due to the overlay was based on preliminary results of a weld overlay optimization study sponsored by the Electric Power Research Institute (EPRI). The residual stress due to a weld overlay depends on the size of the overlay and on whether the direction of interest is hoop or axial. Figure 5.4 shows the hoop direction (axial cracks) residual stress for the worst case without an overlay and for a 1/4 thickness overlay plus the worst case original residual stress.

The IGSCC crack growth rate as a function of applied stress intensity factor is shown in Figure 5.5. The upper bound crack growth law of Figure 5.5 was used for all analyses:

1 l

ga = 4.116 x 10 -12 g 4.615 dT GPC-04-104 20 Revision 1 nutggb

where:

da = differential crack size dT = differential time K = applied stress intensity factor Crack growth as a function of time was calculated by conservatively assuming an infinitely long crack and using the NUTECH computer code NUTCRAK (Reference 8).

The results are shown in Figure 5.6 for an initial crack 0.75 depth of 0.75 inch ( a/t =

1.24

= 0.60 ). Figure 5.6 shows that a 0.75 inch deep crack will grow to a depth of approximately 0.85 inch in five years when the beneficial residual stress due to the weld overlay is considered.

The length of axial cracks is limited by the width of the original butt weld heat affected zone. The weld overlay technique is designed to minimize additional j sensitization by using low weld heat input during the first two layers of weld. Thus the potential for

( additional crack growth in the axial direction is minimized. The maximum axial growth of axial cracks underneath a weld overlay was determined in Reference 13 GPC-04-104 21 Revision 1 nutagh

to be less than 0.01 inch in five years. This axial growth is judged to be insignificant.

.The worst case for an end cap overlay occurs for the end caps that have a crack completely through the original pipe. The crack will not propagate into the overlay weld material due to IGSCC but will grow approximately 0.01 inch due to fatigue in 5 years. Thus, the worst

. case axial crack depth is 1.01 inch which is 79 percent of as-built total wall thickness. Although the crack depth after 5 years exceeds the Reference 2 allowable depth by 4 percent of wall thickness, the weld overlay l design is judged to be acceptable for the following reason. The allowable crack depths in Reference 2 were not allowed to exceed 75% of the wall thickness even though net section collapse analysis would permit much larger cracks for very short crack lengths. This truncation is somewhat artificial and could be eliminated for short, almost through-wall cracks, if leaks are prevented by a weld overlay. Elicination of the truncation results is an allowable depth in excess l

of 79 percent. Thus, the overlay design is acceptable for 5 years.

GPC-04-104 22 Revision 1 nutgsh

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5.1.2.3 Tearing Modulus

' The largest size to which the existing crack could reasonably be expected to grow was postulated to be a one inch radius flaw. This assumes growth of the crack in the radial direction completely through the original

pipe material to the overlay. After such propagation, the assumed crack would be completely surrounded by IGSCC resistant material

the weld between end cap and manifold, the weld overlay, and the end cap and manifold. A tearing modulus evaluation was then performed for this postulated crack. The only applied load was pressure.

The evaluation was performed using the methodology of Reference 11 with material properties from Reference 12.

The postulated flaw and the results are shown in Figure 5.7. The upper dotted line represents the inherent material resistance to unstable fracture in terms of J-integral and Tearing Modulus, T. The line

(

originating at the origin represents the applied loading. Increasing pressure results in applied J-T combination moving up this line, and unstable fracture i

GPC-04-104 23 Revision 1 nutagh

- . . _ - =.

is predicted at the intersection of this applied loading line with the material resistanco line.

Figure 5.7 shows that the predicted burst pressure is in excess of 5500 psig. Thus, there is a safety factor on design pressure of at least 4, which is well in excess of the safety factor inherent in the ASME Code, even in the presence of this worst case assumed crack.

5.2 Elbow Evaluation The largest measured ultrasonic indications in elbow-to-pipe weld number lEll-lRHR-20-BD-3 are an axial crack of depth 94% of wall and length of 3/8 inch, and a circumferential crack of lengen 1-1/2 inches and depth approximately 33% of wall.

5.2.1 Code Stress Analysis 7 A finite element model of the cracked and weld overlaid region was developed using the ANSYS (Reference 6) computer program. Figure 5.8 shows the model. This figure also shows the material that was removed to represent the cracks.

GPC-04-104 24 Revision 1 nutggb

. _. . . . =.-- _, . - . _ . . . - - . .

Based on Reference 4, the applied moments on these welds are:

Weight + OBE Seismic = 743,100 inch-pounds

^

Weight + Steady State Thermal = 636,100 inch-pounds SSE loads are not limiting for the elbow.

j The thermal analysis was performed in the same manner as for the end cap (Section 5.1.1), with appropriate dimensional changes.

The results of a code stress analysis per Reference 1 are given in Table 5.3. The allowable stress values for Reference 1 are also given. The weld overlay repair satisfies the Reference 1 requirements.

A conservative fatigue analysis per Reference 1 was performed. A fatigue strength reduction factor of 5.0

)

was applied due to the crack. The fatigue usage factor i was then calculated assuming 38 startups, 25 small temperature change cycles and one emergency cycle every five years. The results are summarized in Table 5.3.

GPC-04-104 25 Revision 1 i

nutagh

- , , _ ,_ ,......,....,.._n. . - - - - - , . - , , , , , , , . , , , , , , , , , , , , , , ,, , . _ _ , , , -,n.. ,.,_ -,_,

5.2.2 Fracture Mechanics Evaluation Three types of fracture mechanics evaluations were performed. The allowable crack depth was calculated based on Reference 2. Crack growth due to both fatigue and IGSCC was calculated using the NUTECH computer program NUTCRAK (Reference 8) with material constants and methodology from References 9 and 10. Finally, the ultimate margin to failure for a crack assumed to propagate all the way through the original pipe material to the weld overlay was calculated per References 11 and 12.

5.2.2.1 Allowable Crack Depth The allowable depth for a 3/8 inch long axial crack was determined using Reference 2. The dimensions of the repaired elbow were used. Thus, the ratio of applied primary stress to Code allowable stress (S,) was y calculated in the following manner:

l Stress Ratio = g m

GPC-04-104 26 Revision 1 nutagh

P = 1325 psi R = 10.50 inches (Outside Radius of Overlay) i t = 1.16 inches (Overlaid Pipe Thickness)

S, = 16,800 psi Stress Ratio = .71 4

The nondimensional crack length (3) was calculated in .

l the following manner:

i L

- g (rt)1/2 I = .375 inch (Crack Length) g r = 9.87 inches (Mean Radius of Pipe) t = 1.24 inches i = .11 L Thus, per Table IWB-3642-1 of Reference 2, the allowable crack depth is 75 percent of the overlaid wall thickness or a depth of 0.87 inch. Emergency and faulted conditions are not limiting.

GPC-04-104 27 Revision 1 nutgrb

. ~ . .- ._. - __ .- - . - . .. - .. -.

l i

The allowable depth of a 1-1/2 inch long circumferential I crack was also determined using Reference 2. From Table 5.3, the primary stress at the crack location is 16,200 psi. Thus the stress ratio was calculated in the following manner:

16,200 Stress Ratio = Pmg+ Pb = 16,800 = .96 m ,

L l

l The nondimensional crack length was calculated in the following manner: ,

f 1.5 A* * = .02 2 nR 2w(10.5)

Thus based on Table IWB-3641-1 of Reference 2, the j allowable crack depth is 75 percent of the wall l

l thickness. Emergency and faulted conditions are not j limiting.

i l

,i l

t GPC-04-104 28 Revision 1 nutggb

5.2.2.2 Crack Growth i

Crack growth was calculated in a manner similar to Section 5.1.2.2, except: 1) the residual stress due to the weld overlay was changed to represent a one-half thickness overlay; 2) the axial residual stress was used for the circumferential crack.

The axial crack which is essentially through-wall will grow into the IGSCC resistant weld overlay only due to fatigue. The fatigue crack growth for five years of the thermal cycles shown in Figure 5.3 is less than 0.01 inch. Thus, the axial crack depth after five years would be 0.77 inch which is 66 percent of overlaid wall thickness, which is less than the allowable of 75 percent.

Based on Reference 13, the axial growth of the axial crack will be less than 0.01 inch in five years, f

The circumferential crack will grow due to both fatigue and IGSCC. The fatigue crack growth due to five years of the cycles shown on Figure 5.3 is less than 0.01 inch. The IGSCC crack growth was calculated using the upper bound growth curve shown in Figure 5.5 and the GPC-04-104 29 Revision 1 nutggb

residual stress curve shown in Figure 5.9. Crack depth as a function of time is shown in Figure 5.10. Thus, the circumferential crack depth after five years is approximately 0.26 inch which is 23 percent of the overlaid wall thickness which is less than the allowable of 75 percent. Thus, both the worst case axial crack and the worst case circumferential crack will not grow to an unacceptable size within the next 5 years.

5.2.2.3 Tearing Modulus The largest size to which the existing axial crack could reasonably be expected to grow was postulated to be a 0.80 inch radius flaw. This assumes growth of the crack in the radial direction completely through the original pipe material to the overlay. After such propagation, the assumed crack would be completely surrounded by IGSCC resistant material: the weld between elbow and pipe, the weld overlay, a..a the elbow and pipe. A

[

tearing modulus evaluation was then performed for this postulated crack. The normal operating loads of pressure, weight and thermal expansion were applied.

The evaluation was performed using the methodology of Reference 11 with material properties from Reference 12.

GPC-04-104 30 Revision 1 nutagh

The postulated flaw and the results are shown in Figure 5.11. The upper dotted line represents the inherent material resistance to unstable fracture in terms of J-integral and Tearing Modulus, T. The line originating at the origin represents the applied loading.

Increasing load results in applied J-T combinations moving up this line, and unstable fracture is predicted at the intersection of this applied loading line with the material resistance line.

Figure 5.11 shows that the predicted failure load is in excess of 3 times the normal operating loads. Thus, there is a safety factor on normal operating loads of at least 3, which is in excess of the safety factor inherent in the ASME Code, even in the presence of this worst case assumed crack.

5.3 Pipe-to-Pipe Evaluation i

The pipe-to-pipe weld number-lEll-lRHR-24-BR-13 was L determined by ultrasonic examination to have axial crack indications. The largest axial crack is approximately one-half inch long with a depth of approximately 47% of the wall thickness.

GPC-04-104 31 Revision 1

j 5.3.1 Code Stress Analysis

! The weld overlaid regions were assumed to be axisym-metric. That is, a 47% through-wall axial crack was conservatively assumed to be 360 degrees around the pipe and 1/2 inch long centered on the weld. Thus, the assumed crack geometry conservatively envelopes all observed cracks in the pipe-to-pipe weld. In addition, all analyses were conservatively performed using a weld overlay thickness of 0.30 inch which is 25 percent

smaller than the actual thickness of 0.375 inch. A

finite element model of the cracked and weld overlaid region was developed using the ANSYS (Reference 6) computer program. Figure 5.12 shows the model.

Based on Reference 4, the applied thermal, weight and seismic moments on this weld are:

Weight + OBE Seismic = 1,113,000 inch-pounds

(

Weight + Steady State Thermal = 1,626,000 inch-pounds l

SSE is not limiting for this weld. The thermal analysis j was performed in the same manner as for the end cap 1

GPC-04-104 32 Revision 1

_ . . =. _ . _ . . ,

(Section 5.1.1), with appropriate dimensional changes.

The results of a code stress analysis per Reference 1 are given in Table 5.4. The allowable stress values for Reference 1 are also given. The weld overlay repair satisfies the Reference 1 requirements.

A conservative fatigue analysis per Reference 1 was performed. An additional fatigue strength reduction factor of 5.0 was applied due to the crack. The fatigue usage f actor was then calculated with the thermal transients shown in Figure 5.3. The results are

> summarized in Table 5.4, 5.3.2 Fracture Mechanics Evaluation Three types of fracture mechanics evaluations were performed. The allowable crack depth was calculated based on Reference 2. Crack growth due to both fatigue and IGSCC was calculated using the NUTECH computer program NUTCRAK (Reference 8) with material constants and methodology from References 9 and 10. Finally, the ultimate margin to failure for a crack assumed to propagate all the way through the original pipe material to the weld overlay was calculated per References 11 and 12.

GPC-04-104 33 Revision 1 nutagh

5.3.2.1 Allowable Crack Depth The allowable depth for a 1/2 inch long axial crack was determined using Reference 2. The dimensions of the repaired pipe were used. Thus, the ratio of applied primary stress to Code allowable stress (S,) was calculated in the following manner:

Stress ratio =

PR/t g

m P = 1325 psi (Design Pressure)

R = 12.30 inches (Outside Radius of Overlay) t = 1.44 inches (Overlaid Pipe Thickness)

S m

= 16,800 psi (Reference 1)

Substitution yields:

Stress ratio = .67 The nondimensional crack length ('I) was calculated in the following manner:

GPC-04-104 34 Revision 1 nutp_gh

Lg

• • I (rt)1/2 I

f

= 1/2 inch r = 11.58 inches (Mean Radius of Pipe) t = 1.44 inches Substitution yields:

Nondimensional Length = .12 i

I Thus per Table IWB-3642-1 of Reference 2, the allowable crack depth is 75 percent of the wall thickness.

Emergency and faulted conditions are not limiting.

5.3.2.2 Crack Growth Crack growth was calculated in a manner similar to Section 5.1.2.2. The fatigue crack growth for five years of the cycle shown in Figure 5.3 is less than 0.01

! inch. The IGSCC crack growth calculated with the upper bound growth law and an infinite crack length is shown in Figure 5.13. Thus, the axial cracks in the pipe-to-GPC-04-104 35 Revision 1 nutagh

pipe weld will not grow to an unacceptable size ir. the next 5 years.

Based on Reference 13, the axial growth ci the axial crack will be less than 0.01 inch in five years. j 5.3.2.3 Tearing Modulus The largest size to which the existing crack could reasonably be expected to grow was postulated to be a 1.14 inch radius flaw. This assumes growth of the crack in the radial direction completely through the original pipe material to the overlay. After such propagation, the assumed crack would be completely surrounded by i ,

IGSCC resistant material: the pipe-to-pipe weld, the weld overlay, and the annealed piping. A tearing modulus evaluation was then performed for this postulated crack. The applied loads were pressure, seismic, steady state thermal and weight.

The evaluation was performed using the methodology of Reference 11 with material properties from Reference 12.

The postulated flaw and the results are shown in Figure 5.14. The upper dotted line represents the inherent GPC-04-104 36 Revision 1 nutagh

material resistance to unstable fracture in terms of J-integral and Tearing Modulus, T. The line originating at the origin represents the applied loading.

Increasing load results in applied J-T combinations moving up this line, and unstable fracture is predicted at the intersection of this applied loading line with the material resistance line.

Figure 5.14 shows that the predicted failure load is in excess of 4 times the normal loads. Thus, there is a safety factor on normal operating loads of at least 4, which is well in excess of the safety factor inherent in the ASME Code, even in the presence of this worst case assumed crack.

5.4 Sweepolet Evaluation seven small ultrasonic indications were found in the weld between a sweepolet and the 22 inch manifold (weld

( number 1B31-lRC-22AM-1BC-1). All the indications are l transverse to the weld. The largest indication is approximately 1/2 inch long with a depth of approximately 12% of the wall. Figure 5.15 shows the approximate location of the indications.

GPC-04-104 37 Revision 1 nutagh

5.4.1 Code Stress Analysis l

Due to the three dimensional geometry of the sweepolet and the difficulty of performing a repair, a three-dimensional finite element model was developed using ANSYS (Reference 6). The geometry is symmetric about two perpendicular axes. Thus the sweepolet was represented with a 90* model as shown on Figures 5.16 and 5.17. The applied loads are not all symmetric about i both axes. However, the majority of the stress is due to internal pressure which is symmetric about both axes. The applied moments were analyzed by using appropriate symmetric, anti-symmetric or free boundary conditions to represent the full structure of the sweepolet. The stress values presented herein are the highest values at the crack locations.

Based on Reference 4, the applied moments are:

Weight + Seismic = 176,000 inch-pounds Weight + Steady State Thermal = 246,000 inch-pounds GPC-04-104 38 Revision 1 nutagh

The maximum primary stress intensity in the sweepolet is 16,600 psi which is significantly less than the allowable of 25,200 psi.

1 5.4.2 Fracture Mechanics Evaluation Three types of fracture mechanics evaluations were performed. The allowable crack depth was calculated based on Reference 2. Crack growth due to both fatigue and IGSCC was calculated using the NUTECH computer program NUTCRAK (Reference 8) with material constants and methodology from References 9 and 10. Finally, the ultimate margin to failure for a crack of the depth equal to the upper bound predicted depth after an eighteen month fuel cycle was calculated per References 11 and 12.

5.4.2.1 Allowable Crack Depth Due to the three-dimensional state of stress that exists at the sweepolet, the allowable depth was calculated in the same manner as for a circumferential crack.

GPC-04-104 39 Revision 1

Pm + Pb Stress Ratto = S m

Pm + Pb = 16,600 psi (Section 5.4.1)

Sm = 16,800 psi Stress Ratio = 0.99 The nondimensional crack length was calculated in the following manner:

L

- g A "

(rt)1/2 I = .50 inch g

r = 11.0 inches t = .975 inch Thus per Table IWB-3641-1 (Reference 2) the allowable crack depth is 75 percent of the wall thickness.

5.4.2.2 Crack Growth The existing cracks could grow due to both fatigue and stress corrosion. Fatigue growth due to the three types of thermal transients defined in Section 4.2 was GPC-04-104 40 Revision 1 nutagh

. - . - - - _ - - _ _ - - .-~ . . - - - ..

l calculated using the material properties from 4

Reference 9. The fatigue crack growth for five years of the cycles shown in Figure 5.3 was calculated to be less than 0.01 inch.

IGSCC crack growth was calculated using the upper bound j

! crack growth law shown in Figure 5.5. The residual

-l stress distribution normal to the crack is unknown. It was judged that the sweepolet weld residual stress would be equal or less than that due to a butt weld.

Therefore, the residual stress was conservatively assumed to be 30 ksi through-wall bending with tension on the inside surface. The normal stress perpendicular to the crack was determined from the finite element model.

The crack growth analysis was performed per Appendix A of Reference 5. All of the observed cracks are oriented transverse to the sweepolet-to-manifold weld, i

Therefore, the IGSCC crack length is limited to the l

^

width of the heat-affected zone. A finite sized flaw of I

constant length equal to 1/2 inch was assumed. The predicted crack depth as a function of time is shown in Figure 5.18. Maximum crack depth after 5 years is ,

predicted to be 0.38 inch (38 percent of wall l

GPC-04-104 41 Revision 1

. , - - . . - - - c . , - - - - - - - ,,n, 4-, - - - -

-a L

~\

s

\

~.

A N

thickness), which is'well below the allowable of 75 s

.- s percent of wall thickness.

\

5.4.2.3 Tearing Modulus i

t- i <

i %  %.

The largest size to whicn the existing sweepolet crack could reasonably be expected to grow to within one fuel cycle was postulated to be a 0.50 inch radius flaw.

This assumes growth of the crack at a faster rate than the upper bound prediction in Sect' ion 5.4.2.2. A tearing modulus evaluation was then perforned for this postulated crack. The applied loads were pressure, seismic, steady state thernal and weight. ,

\s

'w t The evaluation was performed using the methodology of '

Reference 11 with material properties from Reference 12.

t.

The postulated flaw and the results are shown in Figure 5.19. The upper dotted line represents the inherent ,

material resistance to unstable fracture in terms of ,

~

J-integral and Tearing Modulus, T. The line originating L at the origin represents the applied loading.

Increasing loaoresultsinappipedJ-Tcombinations moving up this line, and unstable fricture is predicted 42 N GPC-04-104 "

Revision 1 w

\

~( nutagh

--a-4 ---,.% e g .--s IM- ,-,, ,- ,-r- - , . , - - -e-i- - - - - - - - - -

s o

at the intersection of this applied loading line with the material resistance line.

Figure 5.19 shows that the predicted failure load is in excess of 3.3 times the normal operating loads. Thus, there is a safety factor on normal operating loads of at least 3.3, which is well in excess of the safety factor inherent in the ASME Code, even in the presence of this worst case assumed crack.

5.5 - Effect on Recirculation and RHR Systems Installation of the weld overlay repairs caused a small amount of radial and axial shrinkage'underneath the

/

overlay. Based on measurements of the weld overlays, the maximum axial shyinkage was 1/4 inch "w (elbow-to-pipe).

,\-

i s *w The' ef fects of the radial shrinkage are limited to the

\ region s

adjacent to and bnderneath the overlay. Based on

\. .

s s.

Reference 13, the stresses due to the radial shrinkage t ,

are Iess than yield stress at distances greater than 4

't m, inches'from the ends of t,he overlay. Weld residual stresses are steady stape secondary stresses and thus (Ire not limited by the ASME Code (Reference 1).

- n 43 GPC-04-104 s ,.

Revision 1 . .

., t . .

nutggh

/

The effect of the axial weld shrinkage on the Recirculation and RHR Systems was evaluated with the NUTECH computer program PISTAR (Reference 14) and the i #

piping model shown in Figure 5.20.

, The four end cap weld overlays are adjacent to free ends of the recirculation manifold. Thus, axial weld shrinkage will not induce stress in any other section of the piping. The measured' axial shrinkage of the elbow weld overlay (.25' inch) and of the pipe-to pipe weld overlay (.19 inch) were imposed as boundary conditions onthismhdel.SincetheASMECodedoesnot limit weld residual stre$s, all c ress indices were set equal to 1.0. ,

i ,

s

't The maximum calculated stress was less than 9 ksi. The location of this stress,.is shown on Figure 5.20. Steady state secondary stresses of 9 ksi are judged to have no

'g ,

I deleterious 'ef fect on the Recirculation or RHR Systems.

~

l t

s s

t f"

'h +.

i Y.

GPC-04-104 44 Revision 1

NORMAL SMALL STARTUP TEMPERATURE EMERGENCY PARAMETER CHANGE CYCLE CYCLE CYCLE (CYCLE 1) (CYCLE 2) (CYCLE 3) 2F U EQUIVALENT 32 F 265 F LINEAR TEMPERATURE AT 1

PEAK 0 8F 640F TEMPERATURE AT 2

THROUGH 368 PSI 8,840 PSI 72,370 PSI WALL THERMAL STRESS e l

Table 5.1 THERMAL STRESS RESULTS GPC-04-104 Revision 1 45

ACTUAL EQUATION STRESS SECTION III CATEGORY NUMBER NB ALLOWABLE OR THICKNESS S N/A N/A S, = 16,800 PSI 10,590 PRIMARY (9) p{

25,200 PSI PRIMARY + 18,950 (10) 50,400 PSI SECONDARY PSI PEAK CYCLE 1 (23,370)5*

CYCLE 2 11) (16,950)5 N/A CYCLE 3 (129,300)5 USAGE FACTOR N/A 0.02 1.0 (5 YR)

• THE FACTOR OF 5 IS THE CONSERVATIVELY ASSUMED FATIGUE STRENGTH REDUCTION FACTOR.

l Table 5.2 END CAP CODE STRESS RESULTS GPC-04-104 Revision 1 46

ACTUAL EQUATION STRESS SECTION III CATEGORY NUMBER NB ALLOWABLE OR THICXNESS S N/A N/A S, = 16,800 PSI 16,200 25,200 PSI PRIMARY (9)

PRIMARY + 19,600 SECONDARY (10) PSI 50'400 PSI PEAK (16,200)S*

CYCLE 1 (8,800)5 (11) N/A CYCLE 2 (83,900)5 CYCLE 3 USAGE FACTOR N/A 0.01 1.0 (5 YR)

• THE FACTOR OF 5 IS THE CONSERVATIVELY ASSUMED FATIGUE STRENGTH REDUCTION FACTOR.

Table 5.3 ELBOW CODE STRESS RESULTS GPC-04-104 Revision 1 47

ACTUAL EQUATION STRESS SECTION III CATEGORY NUMBER NB ALLOWABLE OR THICKNESS S N/A N/A S = 16,800 PSI m

PRIMARY (9) 12,300 PSI 25,200 PSI (10) 16,000 PSI 50,400 PSI C NOARY PEAK CYCLE 1 (19,500)S* N/A (11) (12,950)5 CYCLE 2 CYCLE 3 (125,400)5 USAGE 1 FACTOR N/A 0.019 1.0 (5 YR)

• THE FACTOR OF 5 IS THE CONSERVATIVELY ASSUMED FATIGUE STRENGTH REDUCTION FACTOR.

Table 5.4 PIPE TO-PIPE CODE STRESS RESULTS GPC-04-104 Revision 1 48 nutggh

1 1

M.

MimuM mminM mlumiM M

MlelM MummM MMPE MMM MMM MNM p.mM MMM pgunm MMM MMM4 FPF. l 1

\ll l!.h.

4 m g

--- l hluWm MWW 4MIM Figure 5.1 END CAP FINITE ELEMENT MODEL GPC-04-104 Revision 1 4, nute_c_ h _ - _

- - - - ,_ , - g - - - - _ -

~

l INSULATION 9A

. ,,,,,,,,,,,,,,,,,, , ,,, , ,,,,, 1

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l ', k = 10 BTU /hr-ft- F l

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SECTION A-A Figure 5.2 WELD OVERLAY THERMAL MODEL GPC-04-104 Revision 1 50 00 k

EMERGENCY -

Y SMALL TEMPERATURE -

$ CHANGE STARTUP SHUTDOWN -

NORMAL - 1 J OPERATION l {

RESIDUAL - \ h 38 25 5 CYCLES CYCLES YEARS

u _ +  ; :

~- '

ENTIRE

'1' SEQUENCE CYCLE REPEATS TIME Figure 5.3 THERMAL TRANSIENTS E 0j-104 7

Revision 1 51 g{

40 30 -. PRELIMINARY RESULTS (1/4T) EPRI PLUS WORST CASE WITHOUT

OVERLAY 20 -

10 -

C u

0 , ,

a ID PIPE OD OVERLAY e

cc WORST CASE WITHOUT OVERLAY t

-40 Figure 5.4 AXIAL CRACK GROWTH RESIDUAL STRESS GPC-04-104 Revision 1 52 ggg

10-3 _

6 - P. Ford,1.5 ppm 02 O - R. Horn, 0.2 ppm 02

- 10 4-

. E .

/

da/dT= 1.843 x 10-12K 4.615

_ Upper Bound [

h = 4.116 x 10-12 K4.615 jj/

I 3

'8 I a l O 10  : ,

[ Lower Bound l

_ l l

/

f /

/

10 , , , , ,,,,, , , , ,,,,, , , , , , , , , ,

1 10 20 50 100 1000 Stress intensity Factor (ksi 5)

Figure 5.5 TYPICAL IGSCC CRACK GROWTH DATA (WELD SENSITIZED 304SS IN BWR ENVIRONMENT)

GPC-04-104 Revision 1 53 nutggb

v Alii . . . . .

. .  !!!A

'# # # E4/ # ##

A a

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L  ?

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Niiii :iF

1. 0 - --

y;;; ~ WELDUVERL'IY' O.8 s

0.6 -

a CRACK GROWTH i WITH OVERLAY RESIDUAL STRESS 0.4 -

0. 2 -

l 0.0 , , , .

y 0.0 1.0 2.0 3.0 4.0 5.0 TIME (years)

Figure 5.6 END CAP AXIAL CRACK GROWTH GPC-04-104 Revision 1 54 gg{g

240 -

2CC -

in1b m 150 - JC = 11,200 3 in 2 x 2

^

N 120 -

{ PRESSURE = 6100 PSI b EO -

T

' i JIC = 6000 .2 4g _ \ in PRESSURE = 5500 PSI I -

O . . . . . . .

0 40 80 120 160 200 240 i

OVERLAY WELD

?

/ \

ANNEALED MATERIAL ANNEALED MATERIAL WELD f 1" RADIUS FLAW Figure 5.7 END CAP TEARING MODULUS GPC-04-104 Revision 1 55 nutp_Qh

[ -

'N N 'x '\)'

x TN

sg 's N

x's 's'e

, s s . Ng

, / x\

v N

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z

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Figure 5.8 h

ELBO# # 5 ELENE 06b O

gg,si.S O 56 o@

40 PRELIMINARY RESULTS (1/2 T)

EPRI PLUS WORST CASE WITHOUT OVERLAY 30 -

20 -

10 -

E ID

$ 0 i ,

g OD ELBOW OD g OVERLAY

. G d

E G u

WORST CASE WITHOUT OVERLAY l:

(.

-40 Figure 5.9 CIRCUMFERENTIAL CRACK GROWTH RESIDUAL STRESS GPC-04-104 Revision 1 57 Qd

l l

l 1

r- a n

_/

t

./

1.0 i:N8 WELD OVERLAY 0.8-i:

0.6- ESSENTIALLY NO IGSCC GROWTH PREDICTED u FOR CRACKS OF THIS DEPTH N

m 0.4-PREDICTED IGSCC GROWTH NEXT 5 YEARS 0.2-0.0 . . . .

0.0 1.0 2.0 3.0 4.0 5.0 E(inches)

Figure 5.10 ELBOW CIRCUMFERENTIAL CRACK GROWTH GPC-04-104 58 NU k Revision 1

240 -

200 -

J = 9,700 inib u 150 - in 2 S

x N

• 120 -

3 3 x NORMAL LOADS

=

2.5 x NORMAL

, 80 - LOADS i

40 - UIC = 6000 in 2

0 , , , , , , ,

0 40 80 120 160 200 240 T

OVERLAY WELD

/ \

ANNEALED MATERIAL ANNEALED MATERIAL WELD RdD S FLAW Figure 5.11 ELBOW TEARING MODULUS GPC-04-104 Revision 1 59

1 l

l l

i I

____s (

y ____

t __

Figure 5.12 PIPE-TO-PIPE FINITE ELEMENT MODEL GPC-04-104 Revision 1 60 yy{

y Ai .;u .. .A

' # # # /A// # #/?

L JL

-a Lt

? -

L ,

p ffffffff wi e

1. 0 - - . -

.;;;; ~55I.D6VRLAY' 0.8

h INITIAL DEPTH FOR CRACK TO GROW TO OVEPLAY IN 5 YEARS 0.6 -

Ca 0.4 - PREDICTED IGSCC GROWTH NEXT 5 YEARS

0. 2 -

0.0 , . . ,

i 0.0 1.0 2.0 3.0 4.0 5.0 TIME (years)

Figure 5.13 PIPE TO-PIPE AXIAL CRACK GROWTH GPC-04-104 Revision 1 61 gg

240 -

J = 12,200 i

200 - c .2 in

~ 150 -

1 4 x NORMAL

^

".5 120 -

2

.5 w

s 80 - 3 x NORMAL LOADS "b

40 - J g = 6000 2 0 . , , , , , ,

0 40 80 120 150 200 240 T

< OVERLAY WELD

/ \

ANNEALED MATERIAL ANNEALED MATERIAL WELD / 1.14" RADIUS FLAW Figure 5.14 PIPE-TO. PIPE TEARING MODULUS GPC-04-104 Revision 1 62 nutggh

I v ^ ^v I I I I

[ l

\

c_______ -

u____s N LARGEST INDICATION \

APPROXIMATELY 12% g THROUGH WALL g i

/

i /

h Y D

HOOP , s STRESS /

CRACKS AXIAL Y

STRESS l El VIEW A-A Figure 5.15 SWEEPOLET CRACK GEOMETRY GPC-04-104 Revision 1 63

~

w y N w__ _- #

l %_ l 1

~ - \

~ \\

v \ \

\

/

/ /

, /

/ /

s/ / I s /

i s

Figure 5.16 -

SWEEPOLET FINITE ELEMENT MODEL (OUTSIDE)

GPC-04-104 Revision 1 64 OUk

(

j f - #-

s s

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i

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' I i f f a ' t ! i

', l ,' l l l lll . l l Figure 5.17 SWEEPOLET FINITE ELEMENT MODEL (INSIDE)

GPC-04-104 Revision 1 65 nutech

r* ,r*

// / / / Pb/ / //

T a

1. 1 p m .-

/ / / / ////

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0.4- IGSCC CRACK GROWTH 0.2-i 0.0 . . . .

0.0 1.0 2.0 3.0 4.0 5.0 t

TIME (years)

Figure 5.18 SWEEPOLET CRACK GROWTH GPC-04-104 Revision 1 66

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l ANNEALED MATERIAL ANNEALED MATERIAL 0.5" WELD ] RADIUS FLAW Figure 5.19 SWEEPOLET TEARING MODULUS GPC-04-104 Revision 1 67 nutggh

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Figure 6.20 y PIPING MODEL

6.0 LEAK-BEFORE-BREAK 6.1 Net Section Collapse The simplest way to determine the effect of IGSCC on the structural integrity of piping is through the use of a simple " strength of materials" approach to assess the load carrying capacity of a piping section after the cracked portion has been removed. Studies have shown (References 10 and 12) that this approach gives a conservative, lower-bound estimate of the loads which would cause unstable fracture of the cracked section.

Typical results of such an analysis are indicated in Figure 6.1 (Reference 10). This figure defines the locus of limiting crack depths and lengths for circumferential cracks which are predicted to cause failure by the net section collapse method. Curves are presented for both typical piping system stresses and stress levels equal to ASME Code limits. Note that a very large percentage of pipe wall can be cracked before reaching these limits (40% to 60% of circumference for through-wall cracks, and 65% to 85% of wall thickness for 360* part-through cracks).

GPC-04-104 69 Revision 1 nutg.gh

a. .

s

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, . (

Also shown in Figure 6.1 is'.a samp,1,ing of cracks which

', s have been detected in service, either through UT s

examination or leakage., In each case there has been a comfortable margin between the size crack' that was l t observed and that which-would be predicted to cauSe failure under service loading conditions. Also, as ,

discussed below, there is still considerable margin -

between these net section collapse limits and the actual i,[

cracks which would cause instability.

6.2 Tearing Modulus Analysis Elastic-plastic fracture mechanics analyses are pp presented in Reference 12 which give a more a'ccurat'e representation of the crack tolerance capacity of stainless steel piping than the net section collapse approach described'above. Figures 6.2 and 6.3 c--

graphically depict the results of such an analysis (Reference 12). Through-wall circumferent'lal defects of arc-length equal to 60* through 300* were assumed at various cross sections of a typical BWR Reyirculation System. Loads were applied to these secti'ons of suf ficient magnitude to produce net section,_ limit load, ,

+ N and the resulting values of tearing modulus were

,t compared to that required to cause unstable fracture e v

\ '

\

GPC-04-104 70 -

Revision 1 l

_a nutggh

), ({Qure 6.2). Note that in all. cases there is

-( ~'

substantial margin, indicating ,,that the net section

. s

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collapse limits of the previous s' section are not really

( .

y' , y s

failure limits. Figure 6.3 summarizes the results of r- (

I all such analyses performed for 60* through-wall cracks

^

- in terms of margin on tearing modulus for stability.

The margia. in all cases is substantial.

t+

6.3 Leak Versus Break Flaw Configuration s-4 i, Of pe[ haps more significance to the' leak-before-break s

Ig s s argumcht is the flaw configuration depicted in

,  ; y Figur'e} 6.4. This configurat ion addresses the concerns

' ,\

raised by the occurrence of part-through flaws growing,

- a w Eh respect to the pipe c,ircumference,.before breaking through the outside surface't'o cause leakage. Figure 6.4 presents typical size limitations on such flaws based on the conservative, net section collapse method of Section 6.1. Note that very large crack sizes are predicted. Also shown on this figure are typical

\N l

detectability limits for'short through-wall flaws (which are amenable to leak detecbio{n) and long part-through

- \  : ,

flaws (which dre amenable to Cotection ,v by UT). The n

margins between the detectability limits, and the conservative, net secti'on, collapse failure limits are s

's 3

GPC-04-104 71 Revision 1 D

. nutagh t3

substantial. It is noteworthy that the likelihood of flaws developing which are characterized by the vertical axis shown in Figure 6.4 (full 360' circumferential with no through-wall component) is so remote as to be considered impossible. Material and stress asymmetries always tend to propagate one portion of the crack faster than the bulk of the crack front, which will eventually result in " leak-before-break". This observation is borne out by extensive field experience with BWR IGSCC.

4 6.4 Axial Cracks The recent IGSCC occurrences at Monticello and Hatch 1 were predominately short, axial cracks which grew through the wall but remained very short in the axial direction. This behavior is consistent with expectations for axial IGSCC since the presence of a sensitized weld heat-affected zone is necessary, and this heat-affected zone is limited to approximately 0.25 inch on either size of the weld. Since the major l

loadings in the above net section collapse analysis are bending moments on the cross section due to seismic loadings, and since these loads do not exist in the circumferential direction, the above leak-before-break arguments are even more persuasive for axially oriented l

GPC-04-104 72 Revision 1 nutech

cracks. There is no known mechanism for axial cracks to lengthen before growing through-wall and leaking, and the potential rupture loading on axial cracks is less than that on circumferential cracks.

6.5 Multiple Cracks Recent analyses performed for EPRI (Reference 15) indicate that the occurrence of multiple cracks in a weld, or cracking in multiple welds in a single piping line do not invalidate the leak-before-break arguments discussed above.

6.6 Crack Detection Capability IGSCC in BWR piping is detected through two means: non-destructive examination (NDE) and leakage detection.

Although neither is perfect, the two means complement

, one another well. This detection capability combined with the exceptional inherent toughness of stainless steel, results in essentially 100% probability that IGSCC would be detected before it significantly degraded the structural integrity of a BWR piping system.

GPC-04-104 73 l Revision 1 l

6.7 Non-Destructive Examination The primary means of nondestructive examination for IGSCC in BWR piping is ultrasonics (UT). This method has been the subject of considerable research and development in recent years, and significant improvements in its ability to detect IGSCC have been achieved. Nevertheless, recent In: experience at Monticello, Hatch 1, and else 'ere indicate that there is still considerable room for improvement, especially in the ability to distinguish cracks or crack-like indications from innocuous geometric conditions.

Figure 6.4, however, illustrates a significant aspect of UT detection capability with respect to leak-before-break. The types of cracking most likely to go undetected by UT are relatively short circumferential or axial cracks which are most amenable to detection by leakage. Conversely, as part-through cracks lengthen, and thus become more of a concern with respect to leak-before-break, they become readily detectable by UT, and l

are less likely to be misinterpreted as geometric conditions. This argument is further enhanced by the usual practice of supplementing the UT inspection with radiography (RT) when large UT indications are GPC-04-104 74 Revision 1 nut 9Lh

observed. If a long UT indication is truly a geometric condition, it will be observable as density dif ferences on the radiograph. If, on the other hand, no significant RT density differences are observed in the vicinity of the UT indication, (or if the density differences are abrupt and crack-like), the observed indication is usually diagnosed as IGSCC.

6.8 Leakage Detection Typical leakage detection capability for BWR reactor coolant system piping is through sump level and drywell activity monitoring. These systems have sensitivities on the order of 1.0 gallon per minute (GPM) of unidentified leakage (i.e., not from known sources such as valve packing or pump seals). Plant technical specification limits typically require investigation /

corrective action at 5.0 GPM unidentified leakage.

)

Table 6.1 provides a tabulation of typical flaw sizes to cause 5.0 GPM leakage in various size piping (Reference 10).

Also shown in this table are the critical crack lengths for through-wall cracks based on the net section l

GPC-04-104 75 Revision 1 nutech

collapse method of analysis discussed above. For conservatism, the leakage values are based on pressure stress only, while the critical crack lengths are based on the sum of all combined loads, including seismic.

(Considering other normal operating loads in the leakage analysis would result in higher rates of leakage for a given crack size.) Note that there is considerable margin between the crack length to produce 5.0 GPM leakage and the critical crack length, and that this margin increases with increasing pipe size.

6.9 Historical Experience The above theories regarding crack detectability have been borne out by experience. Indeed, of the approximately 400 IGSCC incidents to date in BWR piping, all have been detected by either UT or leakage, and none have even come close to violating the structural integrity of the piping (Reference 15).

l l

GPC-04-104 76 Revision 1 nutsch

NOMINAL CRACK LENGTH FOR CRITICAL CRACK gg c

PIPE SIZE 5 GPM LCAK (in.) LENGTH 2.e (in.)

4" SCH 80 4.50 6.54 0.688 10" SCH 80 4.86 15.95 0.305 24" SCH 80 4.97 35.79 0.139 I

L Table 6.1 EFFECT OF PIPE SIZE ON THE RATIO OF THE CRACK LENGTH FOR 5 GPM LEAK RATE AND THE CRITICAL CRACK LENGTH (ASSUMED STRESS a = Sm/2)

GPC-04-104 Revision 1 77

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Figure 6.1 l TYPICAL RESULT OF NET SECTICN COLLAPSE ANALYSIS OF l CRACKED STAINLESS STEEL PIPE I GPC-04-104 Revision 1 78 g __-

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Figure 6.4 TYPICAL PIPE CRACK FAILURE LOCUS FOR COMBINED l

THROUGH WALL PLUS 3600 PART THROUGH CRACK GPC-04-104 l Revision 1 81 -

nutach

. - - - - . - . _A.

7.0

SUMMARY

AND CONCLUSIONS The evaluation of the sweepolet flaws and the repairs to the Recirculation and RHR Systems reported herein shows that the resulting stress levels are acceptable for all design conditions. The stress levels have been assessed from the standpoint of load capacity of the components, fatigue, and the resistance to crack growth.

Acceptance criteria for the analyses have been established in Section 3.0 of this report which demonstrate that:

1. There is no loss of design safety margin over that provided by the current Code of Construction for Class 1 piping and pressure vessels (ASME Section III).

2. During the design lifetime of each repair, the observed cracks will not grow to the point where I the above safety margins would be exceeded, f

Analyses have been performed and results are presented which demonstrate that the sweepolet flaws and the GPC-04-104 82 Revision 1 nutp_qh

repaired welds satisfy these criteria by a large margin, and that:

1. The design life of each repair is at least five years.

2. The sweepolet flaws will not grow to an unacceptable size within five years.

Furthermore, it is concluded that the recent IGSCC experienced in the reactor recirculation system at Hatch I does not increase the probability of a design basis pipe rupture at the plant. This conclusion expressly considers the nature of the cracking which has been repaired at Hatch 1, and the likelihood that other similar cracking may have gone undetected. The conclusion is based primarily on the extremely high inherent toughness and ductility of the stainless steel piping material; the tendency of cracks in such piping to grow through-wall and leak before affecting its I structural load carrying capacity (which indeed was the case in the defects observed at Hatch 1); and the fact that as cracks lengthen and are less likely to " leak-before-break", they become more amenable to detection by other NDE techniques such as UT and RT.

1 GPC-04-104 83 Revision 1 nutsch

8.0 REFERENCES

1. ASME Boiler and Pressure Vessel Code Section III, Subsection NB, 1974 Edition with Addenda through Summer 1975.

2. ASME Boiler and Pressure Vessel Code Section XI, Paragraph IWB-3640 (Proposed), " Acceptance Criteria for Austenitic Stainless Steel Piping" (Presented to Section XI Subgroup on Evaluation Standards in November 1982).

3. General Electric Design Specification 22A1344, Revision 3.

4. General Electric letter G-GPC-2-511, "IE Bulletin 79-14 for E. I. Hatch Nuclear Plant Unit 1 -

Transmittal of Preliminary Results of Recirculation System Analysis and Design Drawings,"

December 17, 1982.

5. ASME Boiler and Pressure Vessel Code Section XI, 1980 Edition with Addenda through Winter 1981.

GPC-04-104 84 Revision 1 nutegh

6. ANSYS Computer Program, Swanson Analysis Systems, Revisions 3 and 4. ,

7. Schneider, P. J., " Temperature Response Charts," ..

John Wiley and Sons, 1963.

8. NUTCRAK Computer Program, Revision 0, April 1978, File Number 08.039.0005.

9. EPRI-2423-LD, " Stress Corrosion Cracking of Type 304 Stainless Steel in High Purity Water - a Compilation of Crack Growth Rates," June 1982.

10. EPRI-NP-2472, "The Growth and Stability of Stress Corrosion Cracks in Large-Diameter BWR Piping,"

July 1982.

11. NUREG-0744, Volume 1 for Comment, " Resolution of the Reactor Materials Toughness Safety Issue."

12. EPRI-NP-2261, " Application of Tearing Modulus f Stability Concepts to Nuclear Piping," February 1982. -. . .

l GPC-04-104 85 Revision 1 nutp_qh

• 1

13. NUTECH Report NSP-81-105, Revision 2, " Design Report for Recirculation Line Safe End and Elbow Repair, Monticello Nuclear Generating Plant,"

December 1982.

14. NUTECH Computer Program PISTAR, Version 2.0, Users Manual, Volume 1, TR-76-002, Revision 4, File Number 08.003.0300.

15. Presentation by EPRI and BWR Owners Group to U.S.

Nuclear Regulatory Commission, " Status of BWR IGSCC Development Program," October 15, 1982.

l l

GPC-04-104 86 Revision 1 nutp_qh

9.0 ENGINEERING CHANGES During the installation of the weld overlays, one situation occurred which required an engineering change. The overlay of the pipe-to-pipe weld (IEll-lRHR-24-BR-13) was closer to the pipe to carbon steel valve weld (IEll-lRHR-24-BR-12) than was expected (Figure 9.1). Cracks occurred on the outside surface of the pipe adjacent to the weld overlay (Figure 9.1).

All cracks were removed by grinding to a maximum depth of approximately 0.1 inch. The grinding was extended to approximately the center of the BR-12 weld. The ground out region was then inlaid with NiCrFe weld (Figure 9.2). The weld inlay will produce compressive stresses similar to a weld overlay. Thus the potential for future IGSCC in the stainless steel heat affected zone on the inside of the pipe piece adjacent the BR-12 weld has been significantly reduced by the NiCrFe inlay.

The pipe-to-pipe finite element model was modified to represent the new design (Figure 9.3). The stresses tabulated in Table 5.4 bound the corresponding stresses for the uncracked BR-12 weld.

GPC-04-104 87 Revision 1 nut.e_qh

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